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Influence of interlaminar fracture toughness on impact resistance of glass fibre reinforced polymers
Composites Science and Technology 63 (2003) 943–953 www.elsevier.com/locate/compscitech
Influence of interlaminar fracture toughness on impact resistance of glass fibre reinforced polymers T. Kuboki*, P.-Y.B. Jar, T.W. Forest Department of Mechanical Engineering, University of Alberta, Edmonton, AB, T6G 2G8, Canada Received 23 January 2002; received in revised form 29 August 2002; accepted 17 October 2002
Abstract This work is concerned with delamination resistance of glass fibre reinforced polymers (GFRP) and its influence on GFRP’s resistance to transverse, point impact. The study used two GFRP, one with polyurethane matrix and the other isophthalic polyester. The two GFRPs show distinctly different mode I delamination resistance. Under the same impact condition, at speeds from 1.5 to 5 m/s, the polyurethane-based GFRP developed a smaller damage size than the isophthalic polyester-based counterpart, suggesting that the former has higher impact toughness. However, the two GFRPs showed little difference in the total energy absorbed during the impact, which is another measure commonly used for impact toughness evaluation. We conclude from the study that a consistent trend exists between the delamination resistance in mode I and the critical force for the incipient impact damage. Difference of the impact resistance between the two GFRPs is mainly on the impact damage size developed. The total energy absorbed during the impact remains the same, which is independent of mode I delamination resistance of the GFRP. # 2003 Elsevier Science Ltd. All rights reserved. Keywords: A. Polymer-matrix composites; B. Impact behaviour; B. Fracture toughness; C. Delamination; Polyurethane
1. Introduction Impact resistance is one of the major concerns for laminated fibre-reinforced polymer composites [1–4 as sample references]. The fibre reinforcement provides non-isotropic in-plane strength but produces weak interlaminar resin-rich regions, where under impact loading, extensive damage is generated, especially between plies of different fibre orientation. Impact damage in fibre composites is known to consist of intralaminar matrix cracking, delamination in the interlaminar resin-rich regions, indentation at the contact surface, and fibre breakage, with delamination being the major mode of damage that may cause a significant loss in structural stiffness and lead to catastrophic failure. Therefore, enhancing impact resistance of fibre composites has long been a major task for materials scientists and engineers. In view of the extensive delamination of fibre composites under impact, much work has been dedicated to * Corresponding author. Tel.: +1-780-492-3598; fax: +1-780-4922200. E-mail address: [email protected] (T. Kuboki).
the area. Many studies (for example, [5–10]) have suggested that the delamination under impact is initiated by matrix cracking in a local opening mode. For subsequent growth of a delamination crack, Chang and coworkers [6,7,11] and Springer and co-workers [10,12] suggested that both the opening mode (mode I) and the in-plane shear mode (mode II) are involved. Using a simplified line-loading condition, with (0/90/0) fibre layup to suppress initiation of bending cracks, Razi and Kobayashi [13] analyzed the energy required for the delamination growth in mode II. On the other hand, Sun and Manoharan [5], using (90/0/90) fibre reinforcement, investigated the delamination growth in mode I. Several test methods are widely used to quantify the delamination resistance, among which double cantilever beam (DCB) test is being adopted as an international standard for the measurement of mode I delamination resistance [14]. For the mode II delamination resistance, the 3-point bending is the common mode used to induce fracture, but yet accepted as a standard [15]. Nevertheless, the 3-point bending test on specimens with an end-notched defect, known as end notch flexure (ENF) test, has been used by many researchers to assess composites’ mode II delamination resistance.
0266-3538/03/$ - see front matter # 2003 Elsevier Science Ltd. All rights reserved. doi:10.1016/S0266-3538(02)00316-0
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In contrast to the measurement of delamination resistance, over-all impact resistance of fibre composites has drawn relatively little attention. Early work by Sjo¨blom et al. on carbon fibre reinforced polymers (CFRP) [16] showed that different matrix toughness resulted in difference in the total energy absorbed during the impact, thus suggesting that the impact resistance can be characterized using the total energy absorbed during the impact, provided that the test conditions and test coupon parameters (such as specimen dimensions, fibre volume fraction, etc.) are identical. However, recent work by Jar et al. [17] raised concerns on validity of such an approach to characterize impact resistance of glass fibre reinforced polymers (GFRP). Under the same impact conditions, Jar et al. reported that variation of the matrix toughness did not affect the total energy absorbed during the impact, only causing difference in the damage size development. Therefore, the impact toughness is deemed to vary with the matrix toughness only if the impact toughness is defined as the energy absorbed per unit damage area. This, however, requires a tremendous effort to quantify the damage size and the distribution in the test coupons. As a result, at this point of time it is not clear if such an approach can properly characterize impact resistance of the fibre composites. We recently conducted a series of experiments on mechanical behaviour of GFRP, with an attempt to search for the relationship between the delamination resistance and the impact resistance. This paper compares mechanical behaviour between two GFRP that are widely used in construction industry as artificial timber, and are known to show significantly different resistance to crack development during on-site, construction-related machining, such as sawing and drilling. In this paper, the GFRP’s delamination resistance and impact resistance were characterized using static mode I and mode II delamination tests and drop-weight impact tests, respectively. As to be shown in the following sections, the results elucidate the role of delamination resistance on the GFRP’s impact behaviour.
2. Experimental techniques
provided by ZCL Composites in Edmonton, Canada. The warp unidirectional fabric consists of fibre bundles that are separated by a gap of around 1 mm using stitching threads, to ensure that the bundles are parallel to each other. This type of fabric is a common reinforcement used in oil and gas industries for making GFRP storage tanks. Resin transfer moulding (RTM) technique was used to fabricate the GFRP specimens. The specimens for drop-weight impact tests and those for delamination tests are slightly different in the fibre lay-up, that is [(0/ 90)5]s for the former and [(0/90)402]s for the latter. Both types of the fibre lay-up yield a nominal thickness of 6 mm, but the latter has four central layers with 0 fibre to ensure that the interlaminar crack grows in an environment of uni-directional fibres with the crack growth direction along the fibre orientation. The environment of unidirectional fibre for crack growth, which has been adopted in standards and protocols [18,19], excludes complication in the toughness calculation that may be caused by uncertainty in the crack growth path and the measurement of crack length, thus easing comparison of the material performance. Table 1 summarizes the fibre lay-up and specimen dimensions used in the study. An aluminium insert film of 15 mm thick was placed between 2nd and 3rd of the four central, unidirectional layers to act as the starting defect. During the preliminary study to search for resin injection conditions in the RTM process, we found that the low injection speed adopted in this study did not cause any movement of the insert film. We chose the fibre lay-up of [(0/90)402]s for the delamination test specimens, instead of fully unidirectional fibres, because the fully unidirectional fibre induced a preferential flow of the resin in the RTM process, resulting in panels with a large area of poor resin wetting. The use of (0/90) layers induced isotropic resin flow in the mould to produce specimens of low porosity, as shown in Fig. 1, but the four central layers still provide a unidirectional crack growth environment. Specimens with such a sandwich structure have been successfully used before, to compare delamination resistance of GFRP with a variety of polymer matrix [20]. The previous experience has assured us that
2.1. Materials and specimen fabrication The study used two types of GFRP that are different in the polymer matrix, either isophthalic polyester (TMR300, provided by Viking Plastics, Edmonton) or polyurethane-based resin (PUL-G, provided by Resin System Inc., Edmonton). The isophthalic polyester will be abbreviated as iso-polyester hereafter. The polyurethane-based resin contains 15% CaCO3 particulates that are used to enhance the resin’s stiffness. The glass fibre used was a 9-oz/yd2 warp unidirectional fabric,
Table 1 Fibre lay-up and dimensions for the specimens used in the study Specimen Fibre type lay-up
Specimen dimensions Length of the starting length/width/thickness defect in the specimen (mm) (mm)
93/93/6 [(0/90)5]s Dropweight impact DCB [(0/90)402]s 120/20/6 ENF [(0/90)402]s 120/20/6
–
50 50
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Fig. 1. Optical micrograph of a specimen used in the study. The micrograph was taken from a specimen for delamination test. The aluminium insert film is located in the middle of the left part of the micrograph.
toughness measured using this type of specimens has little difference from that measured from fully unidirectional specimens. Over-all fibre volume fraction of the specimens is 40%, estimated based on the following equation [19]. %Vf ¼
FAW N 100 FD 2h
ð1Þ
where FAW is area weight of the fibre fabric (9 oz/yd2 or 0.3046 kg/m2), N number of fibre layers (equal to 20 in this study), FD fibre density (2560 kg/m3), and 2h specimen thickness (0.006 m). As a nature of the RTM process, the fibre volume fraction was highly reproducible, with variation within 2% among all specimens. 2.2. Mechanical tests The drop-weight impact tests were conducted using Instron Dynatup 8250H instrumented impact tester, equipped with a cylindrical striker of 12.7 mm (1/2 inch) in diameter. Total weight of the tup assembly was 2.69 kg. The specimens were firmly clamped using a pneumatic clamp device, provided by Instron, that has a central circular hole of 76.2 mm in diameter. The impact tests were conducted at speeds of 1.5, 3 and 5 m/s. The results were found to be highly reproducible. Therefore, three specimens were used for each test condition.
The impact tests provided curves of load as a function of time, with signals sampled at a rate of 55 kHz, filtered by an analog filter of 4 kHz. An example of the loadtime curve is shown in Fig. 2(a) that was obtained at an impact speed of 3 m/s. Using the mass of the tup assembly, specimen deflection was calculated based on the Newton’s 2nd Law, as shown in Fig. 2(b). Details of the calculation were given in Refs. [21,22]. The values of load and deflection were then used to produce a loaddeflection curve in Fig. 2(c), or a time function of energy in Fig. 2(d) in which the maximum impact energy and the total absorbed energy were obtained. The total absorbed energy was determined in the following way. Firstly, the deflection point D* where the load was fully released was determined in Fig. 2(c). The corresponding time t* was then determined from the deflection-time curve, Fig. 2(b). With t*, the total absorbed energy was read from Fig. 2(d). The maximum impact energy was read directly from Fig. 2(d). The impact damage was photographed directly under reflected light, as the damage area appeared to be much brighter than the surrounding, which provides good contrast for the photography. The specimen surface that contacts the striker during the impact is named ‘‘front surface’’, and the other side of the specimen is named ‘‘back surface’’. Schematic diagrams for the set-up and the specimen dimensions used for the delamination tests are shown in
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Fig. 2. Illustration of the data obtained from the drop-weight impact test: (a) load vs. time, (b) deflection vs. time, (c) load vs. deflection, and (d) energy vs. time. The curves were from an iso-polyester-based GFRP under an impact speed of 3.0 m/s.
Fig. 3, which adopted most of the specifications provided in ESIS (European Structural Integrity Society) protocols, published in 1993 [19], with some modifications in fibre lay-up, as previously mentioned, and the way in which the piano hinges were bonded to the DCB specimens. The piano hinges were bonded in a way that is opposite to that described in the ESIS protocols, because the RTM mould could only produce specimens of a limited length that if the piano hinges were bonded in the recommended way, a section of around 10 mm would be reduced from the test specimen, resulting in insufficient length for crack growth, estimated to be at least 70 mm for GFRP [23] to establish the resistance curve (R-curve), that is, the plot of GIC versus crack length during the stable crack growth. The choice of piano hinge attachment, however, was found to have little effect on the DCB arm stiffness, and did not affect the calculated toughness value.
Fig. 3. Schematic diagrams of the set-up for (a) DCB test, and (b) ENF test.
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An Instron Universal Testing Machine was used for the DCB and the ENF tests, with a crosshead speed of 1.27 and 2.54 mm/min, respectively. At least 5 specimens were tested for each condition to ensure reproducibility of the results. The DCB test measures the critical strain energy release rate in mode I (GIC) which characterizes the delamination resistance of the GFRP under the opening mode of crack growth. Calculation of GIC was based on the corrected beam theory method [24] in which GIC ¼
3P 2Bða þ jjÞ
E¼
9a2 P 2 16B 2 Eh3
L3 4BCh3
3. Results and discussion 3.1. DCB and ENF tests
ð2Þ
where P is the force, the corresponding displacement, B specimen width, ‘‘a’’ crack length and the x-axis intercept on a plot of C1/3 versus a, which was determined empirically. The GIC values obtained from each DCB specimen are: (i) GIC,init, calculated from the first non-linear point on the force-displacement curve, (ii) GIC,5%, calculated from the point at the 5% off-set from the initial slope on the force-displacement curve, (iii) GIC,prop, defined by the plateau of the R-curve, (iv) GIC,on-set, calculated from the on-set point for unstable crack growth, and (v) GIC,arrest, calculated from the arrest point for the unstable crack growth. The last two GIC values were obtained only for the iso-polyester-based GFRP because its extremely unstable crack growth that resulted in no GIC,5% value and very few points to define the R-curve for GIC,prop. The GIC,on-set and GIC,arrest values offer some indication of delamination resistance of the iso-polyester-based GFRP. Crack growth for the polyurethane-based GFRP, on the other hand, was found to be extremely stable; thus, no GIC,on-set and GIC,arrest values were needed. The mode II ENF test provides the critical strain energy release rate (GIIC) based on the modified beam theory method [25]: GIIC ¼
after the force reached the maximum point, followed by unstable crack growth that caused significant drop of the force. For the ENF test of the iso-polyester-based GFRP, on the other hand, unstable crack growth occurred immediately after the force reached the maximum point.
Results from the DCB test are summarized in Fig. 4 for typical force-displacement curves and Table 2 for GIC values. As shown in Fig. 4, delamination behaviour of the two GFRP was completely different. Crack growth in the polyurethane-based GFRP was very stable, in contrast to the unstable crack growth of the iso-polyester-based GFRP. Table 2 shows that GIC,Init and GIC,Prop for the polyurethane-based GFRP are about 3.7 and 5 times, respectively, of those for the isopolyester-based GFRP, which clearly indicates the former being much tougher in the mode I delamination resistance. A completely different conclusion was drawn from Table 3 that summarizes the ENF test results. The GIIC values suggest that the two GFRP are nearly identical for the mode II delamination resistance. The force-deflection
ð3Þ
ð4Þ
where ‘‘a’’ is the crack length, P the force, B the specimen width, L half the span length, h half the specimen thickness, and 1/C the initial slope of the load-displacement plot. The GIIC values obtained from each ENF specimen are: (i) GIIC,Init, calculated from the first nonlinear point on the force-displacement curve, (ii) GIIC,5%, calculated from the point at the 5% off-set of the initial slope on the force-displacement curve, and (iii) GIIC,Max calculated from the point at the maximum force. It should be noted that, as to be shown in the ‘‘Results and discussion’’, some stable crack growth has occurred in ENF test of polyurethane-based GFRP,
Fig. 4. Typical force-displacement curves of the DCB specimens.
Table 2 Summary of GIC values (in J/m2) and the standard deviation (in the parentheses) from the DCB test Matrix resin Initiation 5% Offset Propagation On-set
Arrest
Isophthalic 436 (55) – 605 (108) polyester Polyurethane 1601 (224) 2352 (310) 3270 (267)
–
609 (110) 309 (65) –
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Table 3 Summary of GIIC values (in J/m2) and the standard deviation (in the parentheses) from the ENF test Matrix
Non-linear
5% Offset
Maximum
Isophthalic polyester Polyurethane
2754 (166) 2766 (628)
3589 (171) 3523 (756)
3530 (168) 3629 (683)
curves from the ENF tests, as shown in Fig. 5, are nearly identical between the two GFRP up to the point of maximum force, after which the iso-polyester-based GFRP showed a sudden force drop, while the polyurethane-based GFRP showed some gradual decrease of the force before a sudden, but less significant force drop. This is an indication of a much more stable crack growth in the polyurethane-based GFRP. However, calculation of the GIIC only used data up to the point of the maximum force. Therefore, GIIC values in Table 3 do not reflect such different fracture behaviour. 3.2. Drop-weight impact test Results from the drop-weight impact tests are summarized in Fig. 6 for typical load-deflection curves, Fig. 7 for typical energy curves, Fig. 8 for the impact damage pattern, and Table 4 for the averaged values of load, deflection, and energy. It should be noted that as shown in Fig. 8, no fibre breakage has occurred in all specimens. All curves in Fig. 6 contain some load oscillation below 4 kN, which is most clearly shown in Fig. 6(a) at the impact speed of 1.5 m/s. Inspection of the postimpact specimens at the speed of 1.5 m/s suggests that little damage, apart from indentation, is visible in polyurethane-based GFRP, as shown in Fig. 8(a). Therefore, the load oscillation below 4 kN is believed to be a perturbation of the GFRP’s response to impact, not
Fig. 5. Typical force-deflection curves of the ENF specimens.
representing the incipient impact damage. This is consistent with that suggested by Knakal and Ireland [26]. At impact speeds of 3.0 and 5.0 m/s, the two GFRP showed significantly different force-deflection curves. For iso-polyester-based GFRP, as shown Fig. 6(b) and (c), a non-recoverable slope change occurred at a load level around 4.5–5 kN. A similar change in slope is also visible in polyurethane-based GFRP, but at a higher load level of around 7–8 kN, as shown in Fig. 6(c). The non-recoverable change of slope in the loaddeflection curves has been suggested by Hirai et al. [27] to represent the incipient damage under impact. They also found the load for the incipient damage to be independent of the impact energy level, and believed that this is a consequence of matrix cracking near the back surface of the specimen [27]. The same point, but on the load-time curve, was identified by Davallo et al. [28] who suggested that the point represents the on-set of delamination, initiated from the matrix cracking. It was noted that in Fig. 6, the load for the incipient damage in the iso-polyester-based GFRP, 4.5–5 kN, happened to be very close to the maximum load applied by the 1.5 m/s impact, and the load for the polyurethane-based GFRP, 7–8 kN, very close to the maximum load applied by the 3.0 m/s impact. Therefore, it is plausible that the slope change in the load-deflection curves due to the occurrence of the incipient impact damage at 1.5 m/s for the iso-polyester-based GFRP and at 3 m/s for the polyurethane-based GFRP exists in Fig. 6, but was not visible because it was too close to the end of the loading section of the curves. Results in Fig. 6 suggest that the critical load for the incipient impact damage in the polyurethane-based GFRP is much higher than that for the iso-polyester-based GFRP. As suggested by Fig. 6(c), difference of the two critical loads is more than 50%. Another major difference between the two GFRP is their impact damage size. Under the same impact condition, the damage size in the iso-polyester-based GFRP is much larger than that in the polyurethane-based GFRP, as shown in Fig. 8. Despite the two GFRP show different critical load for the incipient impact damage, their total absorbed energy, as shown by the curves in Fig. 7, is similar, only differing in the time when the maximum energy is recorded. Jar et al. [17] have observed a similar phenomenon in a previous study, among 4 GFRP of different matrices, in which the absorbed energy was recorded directly using a double-pendulum type impact tester, instead of through the calculation from the force-time curve used in the current study. Consistent results from two independent studies suggest that this phenomenon is worth pursuing further, in order to identify the parameters that control the energy absorption under the impact. This is being carried out in our laboratory.
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Fig. 6. Load-deflection curves of the two GFRP at impact speed of (a) 1.5 m/s, (b) 3.0 m/s and (c) 5.0 m/s. The solid line is from iso-polyester-based GFRP, and the dashed line from polyurethane-based GFRP.
The test results suggest that, based on the criterion of the total energy absorbed during the impact, the two GFRP have similar impact resistance. However, if the criterion also considers the damage area generated during the impact, that is, using the energy absorbed per unit damage area as the criterion, the polyurethanebased GFRP is much tougher than the iso-polyesterbased GFRP. 3.3. Discussion Results from the delamination tests lead to an apparent conclusion. That is, the polyurethane-based GFRP, though with high mode I delamination resistance, does not necessarily show equally high mode II delamination resistance. This conclusion, however, relies on accuracy of the GIIC for characterizing the mode II delamination resistance. As mentioned before, the GIIC values could not reflect the different crack growth behaviour in the
two GFRP, which occurred after the maximum force point was reached. If the crack growth behaviour, such as that shown in Fig. 5, could be quantified, the mode II delamination resistance of the polyurethane-based GFRP would be expected to be higher than the isopolyester-based GFRP. Unfortunately, such a characterization method is yet to be developed. It should be noted that the ENF test did show some difference in GIIC value among CFRP that have either different mode I delamination resistance [29,30] or similar mode I delamination resistance [31]. For GFRP, however, such a trend was not observed [32,33]. Therefore, it is possible that unlike CFRP, GFRP with high mode I delamination resistance may not have equally high mode II delamination resistance. Another possibility is that the current ENF test is not appropriate for the measurement of GIIC, thus suggestions have been made to modify the test method to 4-point ENF test [34]. Whether the current ENF test is appropriate for
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Fig. 7. Energy curve of the two GFRP at impact speed of (a) 1.5 m/s, (b) 3.0 m/s and (c) 5.0 m/s. The solid line is from iso-polyester-based GFRP, and the dot line from polyurethane-based GFRP.
characterizing the GFRP’s mode II delamination resistance or the GFRP has different trend of delamination resistance in mode I and mode II requires further study using other test methods. A possible test method is endloaded split (ELS) test [19] that is known to provide stable crack growth under the mode II loading condition, thus R-curve can be established to characterize the crack growth behaviour. Delamination test results from the current study only allow us to be certain that the mode I delamination resistance of the polyurethane-based GFRP is much higher than the iso-polyester-based GFRP. The impact test results suggest that the total energy absorbed by the specimen is irrelevant to the mode I delamination resistance of the composites, but the load level for the incipient impact damage, that is, the point in Fig. 6(c) where the slope change occurs, is higher for the polyurethane-based GFRP. Since the incipient impact damage has been suggested to be the on-set of delamination from the matrix cracking [28], the results suggest that delamination resistance of the GFRP under
impact is consistent with the mode I delamination resistance measured from the test coupons. The high delamination resistance in mode I for the polyurethanebased GFRP, however, only resulted in the reduction of the damage size, not affecting the total energy absorbed during the impact. The similarity in the total absorbed energy but difference in the impact damage size creates uncertainty in ranking impact toughness of the two GFRP. Using the criterion of the total absorbed energy, the two GFRP show no difference in the impact toughness. But using the criterion of the energy absorbed per unit damage area, the polyurethane-based GFRP becomes tougher than the iso-polyester-based GFRP. The latter criterion, however, requires the measurement of the total damage area. It has been suggested that the total delamination area can be estimated based on the peak impact force value [35–37]. Unfortunately, such a relationship exists only under certain conditions that are not applicable to the two GFRP with different matrix. Therefore, in order
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Fig. 8. Photographs of the impact damages introduced at the speed of (a) 1.5 m/s, (b) 3.0 m/s and (c) 5.0 m/s, in which PI represents the isopolyester-based GFRP and PU the polyurethane-based GFRP.
Table 4 Results from the drop-weight impact test. The standard deviation is shown in the parentheses Impact Matrix resin Max load velocity (kN) (m/s)
Deflection Maximum Total at max impact absorbed load (mm) energy (J) energy (J)
1.5
Polyurethane 4.56 (0.03) 1.47 (0.04) 3.14 (0.07) 1.14 (0.13) Iso-polyester 4.51 (0.03) 1.49 (0.01) 3.11 (0.05) 1.03 (0.05)
3.0
Polyurethane 7.99 (0.06) 2.76 (0.04) 12.40 (0.14) 6.52 (0.07) Iso-polyester 6.73 (0.02) 3.30 (0.07) 12.78 (0.19) 7.25 (0.19)
5.0
Polyurethane 12.51 (0.09) 5.09 (0.19) 35.12 (0.15) 20.51 (0.12) Iso-polyester 11.80 (0.16) 5.72 (0.06) 34.90 (0.29) 20.13 (0.21)
to determine the energy absorbed per unit damage area, the total delamination area needs to be measured directly. Unfortunately, at this stage, there is no established method that assures such a measurement can be made reliably.
There is an additional uncertainty in correlating the impact resistance with the delamination resistance. That is, the DCB and the ENF specimens have crack growth between plies of the same fibre orientation, but the (0/ 90) impact specimens have the crack growth between plies of different fibre orientation. Whether the superior mode I delamination resistance in the unidirectional fibre environment is fully preserved for the crack growth between plies of different fibre orientation needs to be further investigated, in order to clarify the role of the measured delamination resistance from the standard test coupons in the impact resistance of the (0/90) GFRP.
4. Conclusions The study highlights problems that currently exist in the relationship between the delamination resistance of
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GFRP and its impact resistance. The results show a consistent trend between the mode I delamination resistance and the critical force for the incipient impact damage. Difference of the impact resistance between the two GFRP, however, is mainly on the impact damage size developed. The total energy absorbed during the impact remains the same. The results also suggest that despite the difference of the crack growth behaviour in the ENF specimens, the two GFRP show very similar GIIC values. Therefore, another test method should be used to verify whether the difference in the crack growth behaviour is an indication of difference in the mode II delamination resistance between the two GFRP. Acknowledgements The work was sponsored by a research collaboration grant between University of Alberta, Advanced Engineering Materials Centre (Mr. R. W. Fraser) and Resin System Inc., Edmonton (Mr. G. Steadman). We are grateful to the technical support from staff in the Department of Mechanical Engineering, University of Alberta, especially to B. Faulkner for the mechanical testing, A. Yuen for the specimen preparation and T. Hilvo for the photography. References [1] Abrate S. Impact on laminated composite materials. Appl Mech Rev 1991;44(4):155–90. [2] Cantwell WJ, Morton J. The impact resistance of composite materials—a review. Composites 1991;22(5):347–62. [3] Abrate S. Impact on laminate composites: recent advances. Appl Mech Rev 1994;47(11):517–44. [4] Richardson MOW, Wisheart MJ. Review of low-velocity impact properties of composite materials. Composites: Part A 1996;27A: 1123–31. [5] Sun CT, Manoharan MG. Growth of delamination cracks due to bending in a [905/05/905] laminate. Compos Sci Technol 1989;34: 365–77. [6] Choi HY, Wu H-YT, Chang F-K. New approach toward understanding damage mechanisms and mechanics of laminated composites due to low-velocity impact: part II-analysis. J Compos Mater 1991;25:1012–38. [7] Choi HY, Chang F-K. A model for predicting damage in graphite/epoxy laminated composites resulting from low velocity point impact. J Compos Mater 1992;26(14):2134–69. [8] Schoeppner GA, Abrate S. Delamination threshold loads for low velocity impact on composite laminates. Composites: Part A 2000;31:903–15. [9] Karbhari VM, Rydin RW. Impact characterization of RTM composites—II: damage mechanisms and damage evolution in plain weaves. J Mater Sci 1999;34:5641–8. [10] Finn SR, Springer GS. Delaminations in composite plates under transverse static or impact loads—a model. Compos Struct 1993; 23:177–90. [11] Liu S, Kutlu Z, Chang F-K. Matrix cracking and delamination in laminated composite beams subjected to a transverse concentrated line load. J Compos Mater 1993;27(5):436–70.
[12] Finn SR, He Y-F, Springer GS. Delaminations in composite plates under transverse impact loads—experimental results. Compos Struct 1993;23:191–204. [13] Razi H, Kobayashi AS. Delamination in cross-ply laminated composite subjected to low-velocity impact. AIAA Journal 1993; 31(8):1498–502. [14] O’Brien TK. Interlaminar fracture toughness: the long and winding road to standardization. Composites: Part B 1998;29B: 57–62. [15] Davies P, Sims GD, Blackman BRK, Brunner AJ, Kageyama K, Hojo M, et al. Comparison of test configuration for determination of mode II interlaminar fracture toughness results from international collaborative test programme. Plast Rub Compos 1999;28(9):432–7. [16] Sjo¨blom PO, Hartness JT, Cordell TM. On low-velocity impact testing of composite materials. J Compos Mater 1988;22:30–52. [17] Jar P-YB, Gros XE, Takahashi K, Kawabata K, Murai J, Shinagawa Y. Evaluation of delamination resistance of glass fibre reinforced polymers under impact loading. J Adv Mater 2000; 32(3):35–45. [18] American Society for Testing and Materials, ASTM D5528-94a. [19] Protocols for interlaminar fracture testing of composites. Delft (The Netherlands): European Structural Integrity Society (ESIS); 1993. [20] Compston P, Jar P-YB, Davies P. Matrix effect on the static and dynamic interlaminar fracture toughness of glass-fibre marine composites. Composites: Part B 1998;29B:505–16. [21] Wogulis SG, Whitney BW, Ireland DR. Automated data acquisition and analysis for the instrumented impact test. Paper presented at ASTM Symposium on Computer Automation of Materials Testing. Philadelphia, PA. 8 November 1978. [22] Scarponi C, Riotti G, Barboni R, Marcone A, Iannone M. Impact testing on composites laminates and sandwich panels. J Compos Mater 1996;30(17):1873–911. [23] Compston P, Jar P-YB, Davies P. Matrix effect on the static and dynamic interlaminar fracture toughness of glass-fibre marine composites. Composites: Part B 1998;29B:505–16. [24] Hashemi S, Kinloch AJ, Williams JG. Corrections needed in double cantilever beam tests for assessing the interlaminar failure of composites. J Mater Sci Lett 1989;8:125–9. [25] Carlsson LA, Gillespie JW. Mode II interlaminar fracture of composites. In: Friedrich K, editor. Application of fracture mechanics to composite materials. Amsterdam: Elsevier; 1989. p. 113–57. [26] Knakal CW, Ireland DR. Instrumented dart impact evaluation of some automotive plastics and composites. In: Kessler SL, Adams GC, Driscoll SB, Ireland DR, editors. Instrumented impact testing of plastics and composite materials, ASTM STP 936. Philadelphia: American Society for Testing and Materials; 1987. p. 44– 57. [27] Hirai Y, Hamada H, Kim J-K. Impact response of woven glass fabric composites—I. Effect of fibre surface treatment. Compos Sci Technol 1998;58:91–104. [28] Davallo M, Clemens ML, Taylor H, Wilkinson AN. Low energy impact behaviour of polyester-glass composites formed by resin transfer moulding. Plast, Rub Compos Proc Appl 1998;27(8): 384–91. [29] Friedrich K, Walter R, Carlsson LA, Smiley AJ, Gillespie JW. Mechanisms for rate effects on interlaminar fracture toughness of carbon/epoxy and carbon/PEEL composites. J Mater Sci 1989; 24:3387–98. [30] Jar P-Y, Mulone R, Davies P, Kausch H-H. The study of the effect of forming temperature on the mechanical behaviour of carbon fibre/PEEK composites. Compos Sci Technol 1993;46:7–19. [31] Master J E. Correlation of impact and delamination resistance in interleafed laminates. In: Matthews FL, Buskell NCR, Hodgkinson JM, Morton J, editors. Proceedings of 6th International
T. Kuboki et al. / Composites Science and Technology 63 (2003) 943–953 Conference on Composite Materials and Second European Conference on Composite Materials (ICCM-6 & ECCM-2) 1987. Vol. 3:3.96-3.107. [32] Compston P, Jar P-YB, Burchill PJ, Takahashi K. The effect of matrix toughness and loading rate on the mode I interlaminar fracture toughness of glass-fibre/vinyl ester composites. In: Williams JG, Pavan A, editors. ESIS STP27, 2000. p. 37– 48. [33] Compston P, Jar P-YB, Burchill P, Takahashi K. The effect of matrix toughness and loading rate on the mode II interlaminar fracture toughness of glass-fibre/vinyl ester composites. Compos Sci Technol 2001;61(ER2):321–33.
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[34] Davies P, Blackman BR, Brunner AJ. Standard test method for delamination resistance of composite materials: current status. Appl Compos Mater 1998;5:345–64. [35] Jih CJ, Sun CT. Prediction of delamination in composite laminates subjected to low velocity impact. J Compos Mater 1993; 27(7):684–701. [36] Wu E, Shyu K. Response of composite laminates to contact loads and relationship to low-velocity impact. J Compos Mater 1993; 27(15):1443–64. [37] Doxsee LE, Rubbrecht P, Li L, Verpoest I, Scholle M. Delamination growth in composite plates subjected to transverse loads. J Compos Mater 1993;27(8):764–81.
Influence of interlaminar fracture toughness on impact resistance of glass fibre reinforced polymers T. Kuboki*, P.-Y.B. Jar, T.W. Forest Department of Mechanical Engineering, University of Alberta, Edmonton, AB, T6G 2G8, Canada Received 23 January 2002; received in revised form 29 August 2002; accepted 17 October 2002
Abstract This work is concerned with delamination resistance of glass fibre reinforced polymers (GFRP) and its influence on GFRP’s resistance to transverse, point impact. The study used two GFRP, one with polyurethane matrix and the other isophthalic polyester. The two GFRPs show distinctly different mode I delamination resistance. Under the same impact condition, at speeds from 1.5 to 5 m/s, the polyurethane-based GFRP developed a smaller damage size than the isophthalic polyester-based counterpart, suggesting that the former has higher impact toughness. However, the two GFRPs showed little difference in the total energy absorbed during the impact, which is another measure commonly used for impact toughness evaluation. We conclude from the study that a consistent trend exists between the delamination resistance in mode I and the critical force for the incipient impact damage. Difference of the impact resistance between the two GFRPs is mainly on the impact damage size developed. The total energy absorbed during the impact remains the same, which is independent of mode I delamination resistance of the GFRP. # 2003 Elsevier Science Ltd. All rights reserved. Keywords: A. Polymer-matrix composites; B. Impact behaviour; B. Fracture toughness; C. Delamination; Polyurethane
1. Introduction Impact resistance is one of the major concerns for laminated fibre-reinforced polymer composites [1–4 as sample references]. The fibre reinforcement provides non-isotropic in-plane strength but produces weak interlaminar resin-rich regions, where under impact loading, extensive damage is generated, especially between plies of different fibre orientation. Impact damage in fibre composites is known to consist of intralaminar matrix cracking, delamination in the interlaminar resin-rich regions, indentation at the contact surface, and fibre breakage, with delamination being the major mode of damage that may cause a significant loss in structural stiffness and lead to catastrophic failure. Therefore, enhancing impact resistance of fibre composites has long been a major task for materials scientists and engineers. In view of the extensive delamination of fibre composites under impact, much work has been dedicated to * Corresponding author. Tel.: +1-780-492-3598; fax: +1-780-4922200. E-mail address: [email protected] (T. Kuboki).
the area. Many studies (for example, [5–10]) have suggested that the delamination under impact is initiated by matrix cracking in a local opening mode. For subsequent growth of a delamination crack, Chang and coworkers [6,7,11] and Springer and co-workers [10,12] suggested that both the opening mode (mode I) and the in-plane shear mode (mode II) are involved. Using a simplified line-loading condition, with (0/90/0) fibre layup to suppress initiation of bending cracks, Razi and Kobayashi [13] analyzed the energy required for the delamination growth in mode II. On the other hand, Sun and Manoharan [5], using (90/0/90) fibre reinforcement, investigated the delamination growth in mode I. Several test methods are widely used to quantify the delamination resistance, among which double cantilever beam (DCB) test is being adopted as an international standard for the measurement of mode I delamination resistance [14]. For the mode II delamination resistance, the 3-point bending is the common mode used to induce fracture, but yet accepted as a standard [15]. Nevertheless, the 3-point bending test on specimens with an end-notched defect, known as end notch flexure (ENF) test, has been used by many researchers to assess composites’ mode II delamination resistance.
0266-3538/03/$ - see front matter # 2003 Elsevier Science Ltd. All rights reserved. doi:10.1016/S0266-3538(02)00316-0
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In contrast to the measurement of delamination resistance, over-all impact resistance of fibre composites has drawn relatively little attention. Early work by Sjo¨blom et al. on carbon fibre reinforced polymers (CFRP) [16] showed that different matrix toughness resulted in difference in the total energy absorbed during the impact, thus suggesting that the impact resistance can be characterized using the total energy absorbed during the impact, provided that the test conditions and test coupon parameters (such as specimen dimensions, fibre volume fraction, etc.) are identical. However, recent work by Jar et al. [17] raised concerns on validity of such an approach to characterize impact resistance of glass fibre reinforced polymers (GFRP). Under the same impact conditions, Jar et al. reported that variation of the matrix toughness did not affect the total energy absorbed during the impact, only causing difference in the damage size development. Therefore, the impact toughness is deemed to vary with the matrix toughness only if the impact toughness is defined as the energy absorbed per unit damage area. This, however, requires a tremendous effort to quantify the damage size and the distribution in the test coupons. As a result, at this point of time it is not clear if such an approach can properly characterize impact resistance of the fibre composites. We recently conducted a series of experiments on mechanical behaviour of GFRP, with an attempt to search for the relationship between the delamination resistance and the impact resistance. This paper compares mechanical behaviour between two GFRP that are widely used in construction industry as artificial timber, and are known to show significantly different resistance to crack development during on-site, construction-related machining, such as sawing and drilling. In this paper, the GFRP’s delamination resistance and impact resistance were characterized using static mode I and mode II delamination tests and drop-weight impact tests, respectively. As to be shown in the following sections, the results elucidate the role of delamination resistance on the GFRP’s impact behaviour.
2. Experimental techniques
provided by ZCL Composites in Edmonton, Canada. The warp unidirectional fabric consists of fibre bundles that are separated by a gap of around 1 mm using stitching threads, to ensure that the bundles are parallel to each other. This type of fabric is a common reinforcement used in oil and gas industries for making GFRP storage tanks. Resin transfer moulding (RTM) technique was used to fabricate the GFRP specimens. The specimens for drop-weight impact tests and those for delamination tests are slightly different in the fibre lay-up, that is [(0/ 90)5]s for the former and [(0/90)402]s for the latter. Both types of the fibre lay-up yield a nominal thickness of 6 mm, but the latter has four central layers with 0 fibre to ensure that the interlaminar crack grows in an environment of uni-directional fibres with the crack growth direction along the fibre orientation. The environment of unidirectional fibre for crack growth, which has been adopted in standards and protocols [18,19], excludes complication in the toughness calculation that may be caused by uncertainty in the crack growth path and the measurement of crack length, thus easing comparison of the material performance. Table 1 summarizes the fibre lay-up and specimen dimensions used in the study. An aluminium insert film of 15 mm thick was placed between 2nd and 3rd of the four central, unidirectional layers to act as the starting defect. During the preliminary study to search for resin injection conditions in the RTM process, we found that the low injection speed adopted in this study did not cause any movement of the insert film. We chose the fibre lay-up of [(0/90)402]s for the delamination test specimens, instead of fully unidirectional fibres, because the fully unidirectional fibre induced a preferential flow of the resin in the RTM process, resulting in panels with a large area of poor resin wetting. The use of (0/90) layers induced isotropic resin flow in the mould to produce specimens of low porosity, as shown in Fig. 1, but the four central layers still provide a unidirectional crack growth environment. Specimens with such a sandwich structure have been successfully used before, to compare delamination resistance of GFRP with a variety of polymer matrix [20]. The previous experience has assured us that
2.1. Materials and specimen fabrication The study used two types of GFRP that are different in the polymer matrix, either isophthalic polyester (TMR300, provided by Viking Plastics, Edmonton) or polyurethane-based resin (PUL-G, provided by Resin System Inc., Edmonton). The isophthalic polyester will be abbreviated as iso-polyester hereafter. The polyurethane-based resin contains 15% CaCO3 particulates that are used to enhance the resin’s stiffness. The glass fibre used was a 9-oz/yd2 warp unidirectional fabric,
Table 1 Fibre lay-up and dimensions for the specimens used in the study Specimen Fibre type lay-up
Specimen dimensions Length of the starting length/width/thickness defect in the specimen (mm) (mm)
93/93/6 [(0/90)5]s Dropweight impact DCB [(0/90)402]s 120/20/6 ENF [(0/90)402]s 120/20/6
–
50 50
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Fig. 1. Optical micrograph of a specimen used in the study. The micrograph was taken from a specimen for delamination test. The aluminium insert film is located in the middle of the left part of the micrograph.
toughness measured using this type of specimens has little difference from that measured from fully unidirectional specimens. Over-all fibre volume fraction of the specimens is 40%, estimated based on the following equation [19]. %Vf ¼
FAW N 100 FD 2h
ð1Þ
where FAW is area weight of the fibre fabric (9 oz/yd2 or 0.3046 kg/m2), N number of fibre layers (equal to 20 in this study), FD fibre density (2560 kg/m3), and 2h specimen thickness (0.006 m). As a nature of the RTM process, the fibre volume fraction was highly reproducible, with variation within 2% among all specimens. 2.2. Mechanical tests The drop-weight impact tests were conducted using Instron Dynatup 8250H instrumented impact tester, equipped with a cylindrical striker of 12.7 mm (1/2 inch) in diameter. Total weight of the tup assembly was 2.69 kg. The specimens were firmly clamped using a pneumatic clamp device, provided by Instron, that has a central circular hole of 76.2 mm in diameter. The impact tests were conducted at speeds of 1.5, 3 and 5 m/s. The results were found to be highly reproducible. Therefore, three specimens were used for each test condition.
The impact tests provided curves of load as a function of time, with signals sampled at a rate of 55 kHz, filtered by an analog filter of 4 kHz. An example of the loadtime curve is shown in Fig. 2(a) that was obtained at an impact speed of 3 m/s. Using the mass of the tup assembly, specimen deflection was calculated based on the Newton’s 2nd Law, as shown in Fig. 2(b). Details of the calculation were given in Refs. [21,22]. The values of load and deflection were then used to produce a loaddeflection curve in Fig. 2(c), or a time function of energy in Fig. 2(d) in which the maximum impact energy and the total absorbed energy were obtained. The total absorbed energy was determined in the following way. Firstly, the deflection point D* where the load was fully released was determined in Fig. 2(c). The corresponding time t* was then determined from the deflection-time curve, Fig. 2(b). With t*, the total absorbed energy was read from Fig. 2(d). The maximum impact energy was read directly from Fig. 2(d). The impact damage was photographed directly under reflected light, as the damage area appeared to be much brighter than the surrounding, which provides good contrast for the photography. The specimen surface that contacts the striker during the impact is named ‘‘front surface’’, and the other side of the specimen is named ‘‘back surface’’. Schematic diagrams for the set-up and the specimen dimensions used for the delamination tests are shown in
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Fig. 2. Illustration of the data obtained from the drop-weight impact test: (a) load vs. time, (b) deflection vs. time, (c) load vs. deflection, and (d) energy vs. time. The curves were from an iso-polyester-based GFRP under an impact speed of 3.0 m/s.
Fig. 3, which adopted most of the specifications provided in ESIS (European Structural Integrity Society) protocols, published in 1993 [19], with some modifications in fibre lay-up, as previously mentioned, and the way in which the piano hinges were bonded to the DCB specimens. The piano hinges were bonded in a way that is opposite to that described in the ESIS protocols, because the RTM mould could only produce specimens of a limited length that if the piano hinges were bonded in the recommended way, a section of around 10 mm would be reduced from the test specimen, resulting in insufficient length for crack growth, estimated to be at least 70 mm for GFRP [23] to establish the resistance curve (R-curve), that is, the plot of GIC versus crack length during the stable crack growth. The choice of piano hinge attachment, however, was found to have little effect on the DCB arm stiffness, and did not affect the calculated toughness value.
Fig. 3. Schematic diagrams of the set-up for (a) DCB test, and (b) ENF test.
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An Instron Universal Testing Machine was used for the DCB and the ENF tests, with a crosshead speed of 1.27 and 2.54 mm/min, respectively. At least 5 specimens were tested for each condition to ensure reproducibility of the results. The DCB test measures the critical strain energy release rate in mode I (GIC) which characterizes the delamination resistance of the GFRP under the opening mode of crack growth. Calculation of GIC was based on the corrected beam theory method [24] in which GIC ¼
3P 2Bða þ jjÞ
E¼
9a2 P 2 16B 2 Eh3
L3 4BCh3
3. Results and discussion 3.1. DCB and ENF tests
ð2Þ
where P is the force, the corresponding displacement, B specimen width, ‘‘a’’ crack length and the x-axis intercept on a plot of C1/3 versus a, which was determined empirically. The GIC values obtained from each DCB specimen are: (i) GIC,init, calculated from the first non-linear point on the force-displacement curve, (ii) GIC,5%, calculated from the point at the 5% off-set from the initial slope on the force-displacement curve, (iii) GIC,prop, defined by the plateau of the R-curve, (iv) GIC,on-set, calculated from the on-set point for unstable crack growth, and (v) GIC,arrest, calculated from the arrest point for the unstable crack growth. The last two GIC values were obtained only for the iso-polyester-based GFRP because its extremely unstable crack growth that resulted in no GIC,5% value and very few points to define the R-curve for GIC,prop. The GIC,on-set and GIC,arrest values offer some indication of delamination resistance of the iso-polyester-based GFRP. Crack growth for the polyurethane-based GFRP, on the other hand, was found to be extremely stable; thus, no GIC,on-set and GIC,arrest values were needed. The mode II ENF test provides the critical strain energy release rate (GIIC) based on the modified beam theory method [25]: GIIC ¼
after the force reached the maximum point, followed by unstable crack growth that caused significant drop of the force. For the ENF test of the iso-polyester-based GFRP, on the other hand, unstable crack growth occurred immediately after the force reached the maximum point.
Results from the DCB test are summarized in Fig. 4 for typical force-displacement curves and Table 2 for GIC values. As shown in Fig. 4, delamination behaviour of the two GFRP was completely different. Crack growth in the polyurethane-based GFRP was very stable, in contrast to the unstable crack growth of the iso-polyester-based GFRP. Table 2 shows that GIC,Init and GIC,Prop for the polyurethane-based GFRP are about 3.7 and 5 times, respectively, of those for the isopolyester-based GFRP, which clearly indicates the former being much tougher in the mode I delamination resistance. A completely different conclusion was drawn from Table 3 that summarizes the ENF test results. The GIIC values suggest that the two GFRP are nearly identical for the mode II delamination resistance. The force-deflection
ð3Þ
ð4Þ
where ‘‘a’’ is the crack length, P the force, B the specimen width, L half the span length, h half the specimen thickness, and 1/C the initial slope of the load-displacement plot. The GIIC values obtained from each ENF specimen are: (i) GIIC,Init, calculated from the first nonlinear point on the force-displacement curve, (ii) GIIC,5%, calculated from the point at the 5% off-set of the initial slope on the force-displacement curve, and (iii) GIIC,Max calculated from the point at the maximum force. It should be noted that, as to be shown in the ‘‘Results and discussion’’, some stable crack growth has occurred in ENF test of polyurethane-based GFRP,
Fig. 4. Typical force-displacement curves of the DCB specimens.
Table 2 Summary of GIC values (in J/m2) and the standard deviation (in the parentheses) from the DCB test Matrix resin Initiation 5% Offset Propagation On-set
Arrest
Isophthalic 436 (55) – 605 (108) polyester Polyurethane 1601 (224) 2352 (310) 3270 (267)
–
609 (110) 309 (65) –
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Table 3 Summary of GIIC values (in J/m2) and the standard deviation (in the parentheses) from the ENF test Matrix
Non-linear
5% Offset
Maximum
Isophthalic polyester Polyurethane
2754 (166) 2766 (628)
3589 (171) 3523 (756)
3530 (168) 3629 (683)
curves from the ENF tests, as shown in Fig. 5, are nearly identical between the two GFRP up to the point of maximum force, after which the iso-polyester-based GFRP showed a sudden force drop, while the polyurethane-based GFRP showed some gradual decrease of the force before a sudden, but less significant force drop. This is an indication of a much more stable crack growth in the polyurethane-based GFRP. However, calculation of the GIIC only used data up to the point of the maximum force. Therefore, GIIC values in Table 3 do not reflect such different fracture behaviour. 3.2. Drop-weight impact test Results from the drop-weight impact tests are summarized in Fig. 6 for typical load-deflection curves, Fig. 7 for typical energy curves, Fig. 8 for the impact damage pattern, and Table 4 for the averaged values of load, deflection, and energy. It should be noted that as shown in Fig. 8, no fibre breakage has occurred in all specimens. All curves in Fig. 6 contain some load oscillation below 4 kN, which is most clearly shown in Fig. 6(a) at the impact speed of 1.5 m/s. Inspection of the postimpact specimens at the speed of 1.5 m/s suggests that little damage, apart from indentation, is visible in polyurethane-based GFRP, as shown in Fig. 8(a). Therefore, the load oscillation below 4 kN is believed to be a perturbation of the GFRP’s response to impact, not
Fig. 5. Typical force-deflection curves of the ENF specimens.
representing the incipient impact damage. This is consistent with that suggested by Knakal and Ireland [26]. At impact speeds of 3.0 and 5.0 m/s, the two GFRP showed significantly different force-deflection curves. For iso-polyester-based GFRP, as shown Fig. 6(b) and (c), a non-recoverable slope change occurred at a load level around 4.5–5 kN. A similar change in slope is also visible in polyurethane-based GFRP, but at a higher load level of around 7–8 kN, as shown in Fig. 6(c). The non-recoverable change of slope in the loaddeflection curves has been suggested by Hirai et al. [27] to represent the incipient damage under impact. They also found the load for the incipient damage to be independent of the impact energy level, and believed that this is a consequence of matrix cracking near the back surface of the specimen [27]. The same point, but on the load-time curve, was identified by Davallo et al. [28] who suggested that the point represents the on-set of delamination, initiated from the matrix cracking. It was noted that in Fig. 6, the load for the incipient damage in the iso-polyester-based GFRP, 4.5–5 kN, happened to be very close to the maximum load applied by the 1.5 m/s impact, and the load for the polyurethane-based GFRP, 7–8 kN, very close to the maximum load applied by the 3.0 m/s impact. Therefore, it is plausible that the slope change in the load-deflection curves due to the occurrence of the incipient impact damage at 1.5 m/s for the iso-polyester-based GFRP and at 3 m/s for the polyurethane-based GFRP exists in Fig. 6, but was not visible because it was too close to the end of the loading section of the curves. Results in Fig. 6 suggest that the critical load for the incipient impact damage in the polyurethane-based GFRP is much higher than that for the iso-polyester-based GFRP. As suggested by Fig. 6(c), difference of the two critical loads is more than 50%. Another major difference between the two GFRP is their impact damage size. Under the same impact condition, the damage size in the iso-polyester-based GFRP is much larger than that in the polyurethane-based GFRP, as shown in Fig. 8. Despite the two GFRP show different critical load for the incipient impact damage, their total absorbed energy, as shown by the curves in Fig. 7, is similar, only differing in the time when the maximum energy is recorded. Jar et al. [17] have observed a similar phenomenon in a previous study, among 4 GFRP of different matrices, in which the absorbed energy was recorded directly using a double-pendulum type impact tester, instead of through the calculation from the force-time curve used in the current study. Consistent results from two independent studies suggest that this phenomenon is worth pursuing further, in order to identify the parameters that control the energy absorption under the impact. This is being carried out in our laboratory.
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Fig. 6. Load-deflection curves of the two GFRP at impact speed of (a) 1.5 m/s, (b) 3.0 m/s and (c) 5.0 m/s. The solid line is from iso-polyester-based GFRP, and the dashed line from polyurethane-based GFRP.
The test results suggest that, based on the criterion of the total energy absorbed during the impact, the two GFRP have similar impact resistance. However, if the criterion also considers the damage area generated during the impact, that is, using the energy absorbed per unit damage area as the criterion, the polyurethanebased GFRP is much tougher than the iso-polyesterbased GFRP. 3.3. Discussion Results from the delamination tests lead to an apparent conclusion. That is, the polyurethane-based GFRP, though with high mode I delamination resistance, does not necessarily show equally high mode II delamination resistance. This conclusion, however, relies on accuracy of the GIIC for characterizing the mode II delamination resistance. As mentioned before, the GIIC values could not reflect the different crack growth behaviour in the
two GFRP, which occurred after the maximum force point was reached. If the crack growth behaviour, such as that shown in Fig. 5, could be quantified, the mode II delamination resistance of the polyurethane-based GFRP would be expected to be higher than the isopolyester-based GFRP. Unfortunately, such a characterization method is yet to be developed. It should be noted that the ENF test did show some difference in GIIC value among CFRP that have either different mode I delamination resistance [29,30] or similar mode I delamination resistance [31]. For GFRP, however, such a trend was not observed [32,33]. Therefore, it is possible that unlike CFRP, GFRP with high mode I delamination resistance may not have equally high mode II delamination resistance. Another possibility is that the current ENF test is not appropriate for the measurement of GIIC, thus suggestions have been made to modify the test method to 4-point ENF test [34]. Whether the current ENF test is appropriate for
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Fig. 7. Energy curve of the two GFRP at impact speed of (a) 1.5 m/s, (b) 3.0 m/s and (c) 5.0 m/s. The solid line is from iso-polyester-based GFRP, and the dot line from polyurethane-based GFRP.
characterizing the GFRP’s mode II delamination resistance or the GFRP has different trend of delamination resistance in mode I and mode II requires further study using other test methods. A possible test method is endloaded split (ELS) test [19] that is known to provide stable crack growth under the mode II loading condition, thus R-curve can be established to characterize the crack growth behaviour. Delamination test results from the current study only allow us to be certain that the mode I delamination resistance of the polyurethane-based GFRP is much higher than the iso-polyester-based GFRP. The impact test results suggest that the total energy absorbed by the specimen is irrelevant to the mode I delamination resistance of the composites, but the load level for the incipient impact damage, that is, the point in Fig. 6(c) where the slope change occurs, is higher for the polyurethane-based GFRP. Since the incipient impact damage has been suggested to be the on-set of delamination from the matrix cracking [28], the results suggest that delamination resistance of the GFRP under
impact is consistent with the mode I delamination resistance measured from the test coupons. The high delamination resistance in mode I for the polyurethanebased GFRP, however, only resulted in the reduction of the damage size, not affecting the total energy absorbed during the impact. The similarity in the total absorbed energy but difference in the impact damage size creates uncertainty in ranking impact toughness of the two GFRP. Using the criterion of the total absorbed energy, the two GFRP show no difference in the impact toughness. But using the criterion of the energy absorbed per unit damage area, the polyurethane-based GFRP becomes tougher than the iso-polyester-based GFRP. The latter criterion, however, requires the measurement of the total damage area. It has been suggested that the total delamination area can be estimated based on the peak impact force value [35–37]. Unfortunately, such a relationship exists only under certain conditions that are not applicable to the two GFRP with different matrix. Therefore, in order
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Fig. 8. Photographs of the impact damages introduced at the speed of (a) 1.5 m/s, (b) 3.0 m/s and (c) 5.0 m/s, in which PI represents the isopolyester-based GFRP and PU the polyurethane-based GFRP.
Table 4 Results from the drop-weight impact test. The standard deviation is shown in the parentheses Impact Matrix resin Max load velocity (kN) (m/s)
Deflection Maximum Total at max impact absorbed load (mm) energy (J) energy (J)
1.5
Polyurethane 4.56 (0.03) 1.47 (0.04) 3.14 (0.07) 1.14 (0.13) Iso-polyester 4.51 (0.03) 1.49 (0.01) 3.11 (0.05) 1.03 (0.05)
3.0
Polyurethane 7.99 (0.06) 2.76 (0.04) 12.40 (0.14) 6.52 (0.07) Iso-polyester 6.73 (0.02) 3.30 (0.07) 12.78 (0.19) 7.25 (0.19)
5.0
Polyurethane 12.51 (0.09) 5.09 (0.19) 35.12 (0.15) 20.51 (0.12) Iso-polyester 11.80 (0.16) 5.72 (0.06) 34.90 (0.29) 20.13 (0.21)
to determine the energy absorbed per unit damage area, the total delamination area needs to be measured directly. Unfortunately, at this stage, there is no established method that assures such a measurement can be made reliably.
There is an additional uncertainty in correlating the impact resistance with the delamination resistance. That is, the DCB and the ENF specimens have crack growth between plies of the same fibre orientation, but the (0/ 90) impact specimens have the crack growth between plies of different fibre orientation. Whether the superior mode I delamination resistance in the unidirectional fibre environment is fully preserved for the crack growth between plies of different fibre orientation needs to be further investigated, in order to clarify the role of the measured delamination resistance from the standard test coupons in the impact resistance of the (0/90) GFRP.
4. Conclusions The study highlights problems that currently exist in the relationship between the delamination resistance of
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GFRP and its impact resistance. The results show a consistent trend between the mode I delamination resistance and the critical force for the incipient impact damage. Difference of the impact resistance between the two GFRP, however, is mainly on the impact damage size developed. The total energy absorbed during the impact remains the same. The results also suggest that despite the difference of the crack growth behaviour in the ENF specimens, the two GFRP show very similar GIIC values. Therefore, another test method should be used to verify whether the difference in the crack growth behaviour is an indication of difference in the mode II delamination resistance between the two GFRP. Acknowledgements The work was sponsored by a research collaboration grant between University of Alberta, Advanced Engineering Materials Centre (Mr. R. W. Fraser) and Resin System Inc., Edmonton (Mr. G. Steadman). We are grateful to the technical support from staff in the Department of Mechanical Engineering, University of Alberta, especially to B. Faulkner for the mechanical testing, A. Yuen for the specimen preparation and T. Hilvo for the photography. References [1] Abrate S. Impact on laminated composite materials. Appl Mech Rev 1991;44(4):155–90. [2] Cantwell WJ, Morton J. The impact resistance of composite materials—a review. Composites 1991;22(5):347–62. [3] Abrate S. Impact on laminate composites: recent advances. Appl Mech Rev 1994;47(11):517–44. [4] Richardson MOW, Wisheart MJ. Review of low-velocity impact properties of composite materials. Composites: Part A 1996;27A: 1123–31. [5] Sun CT, Manoharan MG. Growth of delamination cracks due to bending in a [905/05/905] laminate. Compos Sci Technol 1989;34: 365–77. [6] Choi HY, Wu H-YT, Chang F-K. New approach toward understanding damage mechanisms and mechanics of laminated composites due to low-velocity impact: part II-analysis. J Compos Mater 1991;25:1012–38. [7] Choi HY, Chang F-K. A model for predicting damage in graphite/epoxy laminated composites resulting from low velocity point impact. J Compos Mater 1992;26(14):2134–69. [8] Schoeppner GA, Abrate S. Delamination threshold loads for low velocity impact on composite laminates. Composites: Part A 2000;31:903–15. [9] Karbhari VM, Rydin RW. Impact characterization of RTM composites—II: damage mechanisms and damage evolution in plain weaves. J Mater Sci 1999;34:5641–8. [10] Finn SR, Springer GS. Delaminations in composite plates under transverse static or impact loads—a model. Compos Struct 1993; 23:177–90. [11] Liu S, Kutlu Z, Chang F-K. Matrix cracking and delamination in laminated composite beams subjected to a transverse concentrated line load. J Compos Mater 1993;27(5):436–70.
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